手機(jī)蓋注塑模具設(shè)計(jì)【一模四腔】【說明書+CAD】

手機(jī)蓋注塑模具設(shè)計(jì)【一模四腔】【說明書+CAD】,一模四腔,說明書+CAD,手機(jī)蓋注塑模具設(shè)計(jì)【一模四腔】【說明書+CAD】,手機(jī),注塑,模具設(shè)計(jì),說明書,仿單,cad International Journal of Machine Tools accepted 4 March 2003 Abstract The use of stereolithography (SL) tooling allows plastic parts to be produced by injection moulding in a very short time due to the speed of mould production. One of the supposed advantages of the process is that it provides a low volume of parts that are the same as parts that would be produced by the conventional hard tooling in a fraction of the time and cost. However, this work demonstrates different rates of polymer shrinkage are developed by parts produced by SL and conventional tooling methods. These revelations may counter the greatest advantages of the SL injection moulding tooling process as the parts do not replicate those that would be produced by conventional hard tooling. This work identifies the different shrinkage that occurs in mouldings produced by an SL mould as compared to those produced from an aluminium mould. The experiments utilise two very different types of polymers and two mould geometries, which are processed in the same manner so that the heat transfer characteristics of the moulds are isolated as the only experimental variable. The work demonstrates how the two mould materials exhibit very different rates of expansion due to the temperature profiles experienced during moulding. This expansion must be compensated for to establish the total amount of shrinkage incurred by moulded parts. The compensation is derived by a mathematical approach and by modelling using finite element analysis. Both techniques depend upon knowledge of the thermal conditions during moulding. Knowledge of these thermal conditions are obtained by real-time data acquisition and simulated by FEA modeling. The application of the findings provide knowledge of the complete shrinkage values relating to the mould material and polymer used which would enable the production of geometrically accurate parts. Keywords: Finite element analysis; Plastic injection moulding; Polymer shrinkage; Rapid tooling without external assistance (i.e. cooling by compressed air) the SL mould would take 15 minutes to return to its ambient tempera- ture. 4.2. Initial shrinkage results The initial shrinkage values, calculated from the part/mould measurements, can be viewed in Table 1. However, these figures need further consideration to establish the total shrinkage that occurs. This is explained in the following section. 4.3. Compensation for thermal expansionby calculation Both of the mould materials used expand when heated, albeit in differing amounts. The measurements Table 1 The % size difference of part/mould, including compensation for thermal expansion by calculation & FEA PA66 Part measurement ABS Part measurement Mould type AL bar SL bar AL disc SL disc AL bar SL bar AL disc SL disc % part/mould difference H110021.3272 H110022.611 H110021.2177 H110022.4149 H110020.7524 H110020.6474 H110020.7351 H110020.5611 % part/mould difference H110021.3434 H110022.7308 H110021.2339 H110022.6101 H110020.7687 H110020.7696 H110020.7514 H110020.7599 including compensation for thermal expansion by calculation % part/mould difference H110021.3841 H110022.7447 H110021.2712 H110022.5649 H110020.8097 H110020.7837 H110020.7887 H110020.7139 including compensation for thermal expansion by FEA taken for shrinkage must be compensated for by the amount of cavity expansion to establish the true amount of difference between mould and part measurements. Not compensating for thermal expansion could lead to an underestimation of shrinkage that occurs during moulding. The value obtained for shrinkage had to be corrected by an amount corresponding to the thermal expansion of the mould. The expansion of the mould resulted in an expansion of the cavity, not contraction. This was verified by plac- ing the mould inserts in an oven at a nominal tempera- ture of 50 C for 10 min and then measuring the cavity in the appropriate measurement axis (see Fig. 2). All moulds showed an increase in cavity size, demonstrating an expansion of the cavity. This was authenticated further in later work (section 4.4.5). This correction factor H9004S M to the part shrinkage due the mould expansion, is given by: H9004S M = am(T m H11002T a ) 13 where: am=linear coefficient of expansion of the mould material in 10 H110026 m/m/K T m =maximum mould temperature during injection cycle in CT a =ambient temperature of the mould on the machine in C. The values used were: L50539 am L50539 SL = 59 10 H110026 m/m/K L50539 AL = 23.8 10 H110026 m/m/K 18 T a : L50539 SL = 23.5 C L50539 AL = 23.5 C. An explanation of these values is given in section 3.4. T m : These values were derived from the thermal history profiles of the moulds and the injection parameters. The temperature at the end of the injection cycle was used in the calculations as this was the point at which pressure application to the injected polymer ceased. When press- 883R.A. Harris et al. / International Journal of Machine Tools & Manufacture 43 (2003) 879887 ure application ended, so did the period at which the shrinkage could be influenced. Any further expansion of the mould after this period would be unable to affect the part size. The derivation of the values is illustrated in Fig. 4. The maximum temperature (T m ) at the end of the injection cycle for each mould type were: L50539 SL disc = 57.46 C L50539 SL bar = 44.37 C L50539 AL disc & bar = 30.39 C. The same maximum temperature was experienced by both the aluminium moulds. This maximum temperature was reached before the end of the injection cycle. The temperature in the stereolithography moulds continued to rise for a long period after completion of the injection cycle. Different maximum temperatures were experi- enced in the SL disc & bar moulds due to the different times of the injection cycle required by their different part volume. Therefore, the calculation for thermal expansion for each mould type was as follows: SL disc mould = 59 10 H110026 (57.46H1100223.5) = 2.00364 mm/m = 0.200364% SL bar mould = 59 10 H110026 (44.47H1100223.5) = 1.23133 mm/m = 0.123133% AL disc & bar mould = 23.8 10 -6 (30.39H1100223.5) = 0.16422 mm/m = 0.016422%. The extra amount of shrinkage determined by calcu- lating the expansion of the mould was incorporated into the percentage shrinkage of the measured parts to reveal the compensated total shrinkage. These values are shown in Table 1. However, these calculations were slightly simplified. An assumption was made that the whole mould was at the temperature determined by the data acquisition for tool temperatures and as such equal Fig. 4. Thermal conditions in the moulds during injection cycle. expansion was experienced throughout. In reality the increase in temperature was localised around the mould- ing cavity. Further work was required to prove whether any differing expansion of the mould cavity occurred due to the temperature distribution and to assess any such effects on the compensation values used. 4.4. Compensation for thermal expansionby Finite Element Analysis Finite Element Analysis (FEA) was used to model the mould expansion due to the localised heating caused by the hot polymer contained within the cavity. The FEA software package used in this work was Algor. Two forms of FEA analysis from Algor were used. Firstly, a transient thermal analysis was conducted in order to determine the temperature distribution in the mould. Then a linear elastic analysis was performed using the temperature distribution results to determine the resulting expansion of the mould cavity. 4.4.1. FEA modelling. Stage 1creation of model The FEA work began with the creation of the required models and FEA mesh. In order to reduce solution time, one quarter of the full mould was created, due to the quarterly symmetrical nature of each of the specimens gauge length. This is shown in Fig. 5. The model was extruded such that the mesh spacing between each node was 0.5 mm in the immediate region beneath the moulding cavity (4 mm deep) as these are the elements of interest when evaluating the cavity expansion. This was conducted to ensure that a node was at the equivalent point to the thermocouple in the experi- ments, from which the critical time dependant tempera- tures were derived (see section 3.3). After this the mesh spacing for the rest of the model was 10 mm in order to achieve a shorter analysis time. 4.4.2. FEA modelling. Stage 2allocation of material properties The materials were assumed to be homogeneous and isotropic with constant material properties independent of temperature. The values are listed in Table 2. 4.4.3. FEA modelling. Stage 3transient thermal analysis Transient thermal analysis refers to a thermal con- dition where temperature is a function of time. This analysis type was relevant to the conditions that occurred within the moulds during the experiments. The heat was supplied by the injected polymer which transferred its energy (heat) into the surrounding mould material. This energy was not limitless and the polymer corresponded by reducing in temperature as the heat was transferred into the lower temperature mould. 884 R.A. Harris et al. / International Journal of Machine Tools & Manufacture 43 (2003) 879887 Fig. 5. FEA model section of bar & disc moulds. Table 2 Material values used in FEA for epoxy, aluminium & polymer a Epoxy SI units Reference source r (density) 1250 kg/m 3 22 K (thermal conductivity) 0.19 W/m K 22 C p (specific heat capacity) 1046.7 J/kgC 22 E (modulus of elasticity) 2.610 9 N/m 2 22 n (poissons ratio) 0.35 22 a (thermal coefficient of expansion) 5910 H110026 /C 18 Aluminium r 2720 kg/m 3 23 K 170 W/m o K 23 C p 880 J/kgC 23 E 68.910 9 N/m 2 23 n 0.3 23 a 2310 H110026 /C 23 Polymer r 1145 kg/m 3 22 K 0.2962 W/m K 22 C p 1625.7 J/kgC 22 E 0.2610 9 N/m 2 ( a )( a ) n 0.35 22 a 59 or 2310 H110026 /C( a ) 18 or 23 a The value of E for the polymer was a nominal low value to reduce restraint. The value of a in the polymer was made the same as the mould material to prevent interference. The use of such values in both cases was to ensure that the presence of the polymer did not influence the expansion of the cavity in the simulation To conduct the transient thermal analysis, certain assumptions were made: L50539 The polymer was all initially at 270 C and was in perfect contact with the mould at all times. L50539 There was no thermal resistance between the plastic and the mould. L50539 The mould material was all initially at 23.5 C. The critical time dependant temperatures (see Fig. 4) at the positions equivalent to thermocouple placement in the experiments were plotted from the FEA results and the correct step in the solution noted. From this tempera- ture distribution it was possible to investigate the reac- tion of the model (expansion) to a chosen condition (temperature) at a pre-identified time step by performing a linear elastic analysis. 885R.A. Harris et al. / International Journal of Machine Tools & Manufacture 43 (2003) 879887 4.4.4. FEA modelling. Stage 4linear elastic analysis The models were fully restrained on the planes of symmetry, on the basis of which the FEA models were devised (shown in Fig. 5) and the parting plane of the mould inserts. No further restraints were applied allowing free expansion. The latter point may at first seem a little odd as the mould insert was contained in a steel pocket within the injection moulding bolster. However it has been already shown that even in opti- mum conditions the maximum possible expansion of the whole insert could be just 0.25 mm in any one direc- tion. This amount of expansion could occur freely due to the clearance required to enable fitting and removal of the inserts in the bolster pocket. The planes of restraint are illustrated in Fig. 6. 4.4.5. FEA modelling. Stage 5results After the linear elastic analysis was performed the movements (thermal expansion) of the model rep- resenting the planes of measurement (see Fig. 2) were determined by interrogating the displacement vectors of the nodes representing the cavity edge. An example of this is illustrated in Fig. 7. From these displacements an average movement (expansion) was determined. In each case the results demonstrated an outwards expansion of the cavity. To determine the total mould expansion over the measure- ment axis, the average figure derived from the displace- ment vectors was doubled to include axial expansion in the opposite direction, as only half of the measurement axis was modelled. The results of these expansion values on total part shrinkage are detailed in Table 1. 5. Discussion It can be seen from the results (Table 1) that allowing for thermal compensation reveals higher values of Fig. 6. Planes of restraint on bar & disc FEA models. Fig. 7. Interrogation of displacement vectors in FEA results to estab- lish amount of expansionexample shows AL bar. shrinkage in all polymer/tool combinations, albeit at dif- fering scales. Some of the values revealed quite different shrinkage values when the mould expansion was incor- porated. The FEA values gave a good comparison with those derived from the calculations. This gave confidence that the method used did give an accurate assessment of the mould expansion and the resulting total shrinkage of the parts. It should be noted that the FEA method is an 886 R.A. Harris et al. / International Journal of Machine Tools & Manufacture 43 (2003) 879887 approximate solution and relies on the accuracy of the model, mesh density, material properties, surface contact resistance and physical restraints. Although many of the figures used in the FEA were definitive as they were produced from real practice, some were the result of assumptions that had to be made. When compensation for thermal expansion is included, several themes in the part shrinkage results from the experimental data were revealed. These include: 5.1. Shrinkage direction The crystalline polymer (PA66) demonstrated slightly greater shrinkage differences (7% more) in the polymer flow direction (bar specimen) as opposed to the direction perpendicular to polymer flow (disc specimen). This was also shown with the amorphous (ABS) parts but to a lesser degree (3% more). These characteristics are typi- cal of all injection moulded parts, with crystalline parts being more susceptible to anisotropy (directional differences) due to the flow direction causing greater alignment of chains within the polymer 19. 5.2. Shrinkage according to mould materialPA66 The results show that the shrinkage that occurred in PA66 parts from the SL moulds of both geometries was double that incurred by the comparative parts from the AL moulds. An expected shrinkage range for PA66 is 1 2.2% 20. The parts from the AL moulds demonstrated shrinkage just above the minimum amount expected, while the parts from the SL moulds incurred shrinkage above the maximum in the expected shrinkage range. Also the range of part of sizes measured were much greater than those experienced in the ABS parts. PA66 part measurements differed in a range of 0.35 mm as compared to 0.18 mm for the ABS parts. This is a typi- cal characteristic of crystalline polymers which are more difficult to hold part tolerances, as compared to amorph- ous polymers 21. 5.3. Shrinkage according to mould materialABS The results show that the shrinkage of the ABS parts is largely unaffected by the mould material used. A shrinkage of 0.76% was incurred by all ABS parts in the experiments. An expected shrinkage range for ABS is 0.50.6% 20. 6. Conclusions This work has defined that double the amount of shrinkage occurred in PA66 (a crystalline polymer) when injection moulded in an SL tool, as compared to an AL tool. In the same experimental conditions ABS (an amorphous polymer) demonstrated no such differ- ences. The importance of compensating for thermal expan- sion of the mould in the calculation of shrinkage has been demonstrated. This is critical in determining absol- ute shrinkage values in plastic tools, which expand more than metal tools. Neglecting the mould expansion in plastic tools would lead to significant error in determin- ing the absolute part shrinkage. The establishment of differing part shrinkage in crys- talline polymers exposes a flaw in the use of shrinkage compensation factors supplied by manufacturers. This work has shown that the shrinkage of crystalline poly- mers is dependant upon process conditions which are variable. Supplied shrinkage factors would be specific only to the conditions under which the test pieces were produced. Thus, traditional shrinkage factors are insuf- ficient not only in the use of SL tools, but also any other techniques where there is any significant process vari- ation from the norm. It has been shown that the shrinkage of an amorphous polymer was unaffected by the cooling conditions which were imposed by mould material type. Consequently, where possible, it is recommended that amorphous poly- mers are used in preference to crystalline alternatives when using SL moulds. References 1 P.D. Hilton, P.F. Jacobs, Rapid Tooling: Technologies and Indus- trial Applications, Marcel Dekker, 2000. 2 T. Luck, F. Baumann, U. Baraldi, Comparison of downstream techniques for functional and technical prototypesfast tooling with RP, Proceedings of Fourth European RP Conference, 13 15 June, 1995, Belgriate, Italy, pp. 247260. 3 J. Eschl. Experiences with photopolymer inserts for injection moulding, European Stereolithography Users Group meeting, 2 5 November, Florence, Italy, 1997. 4 R.A. Harris. Direct AIM tooling. Rapid prototyping & tooling state of the industry annual worldwide progress report, Wohlers Report 2002, Published by Wohlers Associates Inc, USA, Part 3: Tooling, p. 70, 2002. 5 A. Schulthess, B. Steinmann, M. Hofmann, Cibatool SL epoxy resins and some new applications. Proceedings of the 1996 North American Stereolithography Users Group Meeting, 1014 March, San Diego, CA, 1996. 6 M. Damle, S. Mehta, R. Malloy, S. McCarthy, Effect of fibre orientation on the mechanical properties of an injection molded part and a stereolithography-insert molded part. Proceedings of the Society of Plastics Engineers (SPE) Annual Technical Confer- ence (ANTEC), Atlanta, GA, pp. 584588, 1998. 7 D.A. Velarde, M.J. Yeagley, Linear shrinkage differences in injection moulded parts, Plastics Engineering, The Society of Plastics Engineers, December 2000, pp. 6064, 2000. 8 P. Gipson, P. Grelle, B. Salamon, The effects of process con- ditions, nominal wall thickness, and flow length on the shrinkage characteristics of injection molded polypropylene, The Journal of Injection Molding Technology 3 (3) (1999) 117125. 9 P. Patel, Effect of processing conditions on the shrinkage and 887R.A. Harris et al. / International Journal of Machine Tools & Manufacture 43 (2003) 879887 crystallinity of injection moulded parts, Proceedings of the Society of Plastics Engineers (SPE) Annual Technical Confer- ence (ANTEC), Toronto, Canada, pp. 36323635, 1997. 10 D. Tursi, S.P. Bistany, Process and tooling factors affecting sink marks for amorphous and crystalline resins, Journal of Injection Molding Technology 4 (3) (2000) 114119. 11 D. Pierick, R. Noller, The effect of processing conditions on shrinkage. Proceedings of the Society of Plastics Engineers (SPE) Annual Technical Conference (ANTEC), Montreal, Canada, pp. 252253, 1991. 12 BS EN ISO 294-1. Plasticsinjection moulding of test speci- mens of thermoplastic materialsPart 1: General principles, and moulding of multipurpose and bar test specimens, British Stan- dards Institution, Issue 2, ISBN 0 580 27299 0, 1998. 13 BS EN ISO 294-4. Plasticsinjection moulding of test speci- mens of thermoplastic materialsPart 4: Determination of moulding shrinkage, British Standards Institution, Issue 2, ISBN 0 580 27826 3, 1998. 14 ASTM D955, Standard test method for measuring shrinkage from mold dimensions of molded plastics, American Society for Test- ing and Materials, 1996. 15 R